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TOUGHNESS IN HY-80 GMA WELDS

T. A. Siewert

R. E. Trevisan

R T. Purtscher

National Bureau of Standards

U.S. Department of Commerce

Boulder, Colorado 80303

April 1986

NBSIR 86-3043

FUSION LINE SHAPE VERSUS

TOUGHNESS IN HY-80 GMA WELDS

T. A. Siewert

R. E. Trevisan

R T. Purtscher

Fracture and Deformation Division

Institute for Materials Science and Engineering

National Bureau of Standards

Boulder, Colorado 80303

April 1986

Supported by

D. W. Taylor Naval Ship R&D Center

Department of the Navy

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige. Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler. Director

CONTENTS

Page

ABSTRACT 1

1. INTRODUCTION 2

2. EXPERIMENTAL DESIGN 3

3. PROCEDURE 4

3.1 Materials 4 3.2 Weld Procedure 4 3.3 Specimens 5 3.4 Tests 6

4. RESULTS AND DISCUSSION 8

5. CONCLUSIONS 12

6. ACKNOWLEDGMENTS 12

7. REFERENCES 13

Figures 15

Tables 34

iii

Fusion Line Shape versus Toughness in HY-80

GMA Welds

T. A. Siewert, R. E. Trevisan, and P. T. Purtscher

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado

This report describes the effect of the electrode weave procedure on

both the fusion-line shape and toughness of reduced-gap gas metal arc (GMA)

welds. To study the variation in toughness, four GMA welds were produced in

25-mm-thick HY-80 plate using MIL 100S-1 electrode and following the weld

procedure listed in the electrode specification, MIL-E-23765/2C . These four

welds were used to compare stringer beads with various weave procedures using

both manually controlled and adaptively controlled welding systems.

All-weld-metal Charpy V-notch (CVN) and tensile tests confirmed that these

welds met the requirements for MIL 100S-1 electrodes. An increase in

strength from 630 to 680 MPa and a decrease in toughness from 135 to 110 J were observed for the weave patterns with the greatest penetration into the

base material.

Testing of compact tension and single-edge-notch bend specimens at room temperature produced Kq(J) fusion-line toughness values of 180 to 315 MPaVm,

with little variation due to weave procedures. Microstructural examination

of the crack-initiation region indicated failures occurred and propagated in

the HAZ or in the weld metal, rather than along the fusion line. An etching

technique revealed a small number of planar flaws on the fusion-line surface

that were not detected with radiography. We conclude that weave procedure

substantially modifies the fusion-line shape, can change the weld strength,

and has little effect on fusion-line or weld toughness.

Key words: electrode weaving; fracture toughness; fusion line; gas metal

arch welding; HY-80; mechanical properties; weld properties.

1
1 . INTRODUCTION

Many factors influence the mechanical properties of welds, including

welding processes, conditions, and techniques. The magnitude of the effect

of changing the welding process, a factor that has been widely studied and

reported, can be estimated by comparing reported Charpy V-notch (CVN)

absorbed energies for HY-80 welds produced by two welding processes. Specification requirements and toughness of welds produced by the GMA and SMA

processes are compared in table 1 [1-9]; base material toughness is included

for reference [10,11],

Weld toughness can be affected substantially by the welding process, as indicated by the factor-of-two difference between GMA and SMA welds. Also,

all the weld toughness values in table 1 are well below those for the base

material indicating that the weld toughness controls the fitness of the

entire welded structure. Therefore, the weld merits special attention in

toughness tests.

Many variables (other than the welding process) affect toughness, but

the effects of some have not been measured. This knowledge is necessary for

production of welds with optimum properties. The weld weave procedure

(especially of reduced-gap welds) is one of these variables for which little

information is available and therefore was the subject of this study.

HY-80 is one of the alloys that can have a very complicated interface

between the weld and heat affected zone (HAZ) [12,13] owing primarily to the

significant compositional differences of the weld metal, the weld metal that

is diluted with molten base metal, and the base metal. Also, as the weld

cools and begins to solidify, the last layer of the base plate to melt may

not have time to mix with the weld metal. This "unmixed zone" has the base

metal composition but is subjected to the weld cooling cycle. This rapid

cooling may also result in very different mechanical properties in the

various weld regions.

Study of the weld's centerline toughness and strength does not reveal the

total effect of weld weave procedure on the toughness. Although such tests

of the centerline properties form an important part of this study, the tough- ness of the region near the HAZ and its effect on the failure mode was also

measured. The toughness of the HAZ is particularly important for reduced-gap

2
welds since the reduced bevel angle reduces the through-thickness dimension

of this "unmixed zone" region, making it an easier fracture path if it has

low toughness.

2. EXPERIMENTAL DESIGN

Welding gun oscillation amplitude and arc penetration into the plate

edges forming the sides of the weld joint were the variables in the study.

We used two levels of oscillation amplitude (none and 10 mm), minimum and

high penetration, and two welding systems to produce the four welds.

Following the experimental design, summarized in table 2, we evaluated a progression of electrode weave techniques, extending from a singledimensional mechanical gun oscillator through a fully automated welding

system that controlled all three axes of gun motion.

The stringer bead and simple gun oscillator are typical of current manufacturing technology. These systems must be continually monitored, since

joint variations due to inaccurate fit or thermal stresses require constant

adjustment of position. The fully automated system produced a more consistent joint through adaptive control of the gun position during each weave

cycle. Weld 1 was made with a motorized carriage that produced a series of

stringer beads. Weld 2 incorporated a 10-mm-wide horizontal weave using the

mechanical gun drive to improve fusion into the sides of the weld joint.

This weld was produced with minimum penetration of the arc into the sides of

the weld joint as was weld 1 . Weld 3 was like weld 2 but used a larger

oscillation-amplitude to increase penetration of the arc into the sides of

the weld joint. This was accomplished by adjusting the horizontal oscillation

so that it went further into the side of the weld joint, simulating the

action of a welding system that senses the sides of the joint through a reduced arc voltage. Weld 4 was produced with such an automated system, one

that followed the weld joint.

3
3. PROCEDURE

3.1. Materials

A single heat of 25-mm-thick HY-80 steel plate conforming to MIL-S-16216J

Amendment 1 was used for the base material. Its chemical analysis, mechanical properties, and heat identification are listed in table 3* The plate was

flame cut into 0.25-m x 0.75-m pieces, all aligned with the plate rolling

direction so that the HAZ texture would not vary among the welds. These

plates were joined in pairs along their 0.75-m sides to form the weld joints.

The flame-cut edges and plate surfaces adjacent to the joint were handground

to remove the surface oxides. A matching composition backing strip, 0.8 m x 25 mm x 5 mm, was used to support the first weld pass.

To facilitate testing of the fusion line, a 45° included-angle singlebevel joint was used. Figure 1 shows the joint configuration and plate

dimensions. The gap between the two plates was 10 mm at the weld root.

Figure 2 shows the plate alignment on the welding table.

The welding gun was angled 20° to one side of vertical to promote good

fusion to both sides of the single-bevel joint and to provide gas nozzle

clearance, as shown in figure 3. A single heat of MIL 100S-1 electrode,

1.2 mm in diameter and conforming to MIL-E-23765/2C was used for all four

welds. Its chemical composition, mechanical properties, and heat identification are listed in table

3.2 Weld Procedure

The plates (weld joints) were preheated to a temperature between 120 and

150°C to prevent hydrogen-induced cracking. The plates were not allowed to

cool below 120°C before the next pass was begun. If a weld was only partially completed in one shift, it was preheated the next morning before the

welding was continued. After the completion of welding, the test plate was allowed to cool in still air, a relatively slow process for the large

( 80-kg-average ) test plates. Welding was performed in the flat ( 1 G) position. Usually, each joint required ten weld passes, and all passes were made

in the same direction. Other welding conditions are summarized in table 5.
The first pass was produced with the lowest possible current and arc

voltage. This produced good fusion at the sides of the joint, but minimized

dilution by the plate material, particularly by the joint backing strip. The

low dilution reduced the amount of carbon and chromium from the plate

material that would mix in the weld; the resulting lower carbon equivalent

ensured formation of a satisfactory low hardness-high toughness microstructure in the first weld pass.

The subsequent passes were performed at a higher current and arc voltage, which produced a more stable and controllable, albeit more penetrating,

arc. The greater penetration was primarily into previously deposited weld

metal and therefore could be tolerated. The travel speed of these subsequent

passes was increased to keep the heat input similar to that of the first pass

(2.7 kJ/mm).

Special attention was given to the quality of the weld on the vertical

side of the joint since this was the side used in fracture testing. The

first weld bead for each layer was produced on the vertical side to minimize

any incomplete fusion defects due to weld metal flowing ahead of the arc. The heat input was kept constant so that the only differences among the welds

were those due to the oscillation pattern. The typical oscillation patterns

used in each of the three welds are summarized in table 6.

3-3. Specimens

Specimens for tensile, CVN , compact tension (CT), and single-edge-notch

bend (SENB) testing were machined from the welded plates, whose orientations

are shown schematically in figure 4. The tensile specimens were prepared

according to ASTM specification E 8. The CVN specimens were prepared according to ASTM E 23, and the notches were oriented perpendicular to the plate

surface and such that the crack would propagate in the direction opposite to

that of welding. The CT and SENB specimens were prepared according to

ASTM E 399, and the notches were oriented perpendicular to the plate surface

and such that the crack would propagate in the direction of welding. In the

following discussion, these specimens are grouped according to whether they were used to evaluate the weld metal (tensile and CVN) or the fusion line

(CT and SENB).

5
The specimens are identified with a three-term name: the first indicates

the weld number, the second indicates the specimen type (e.g., CT, SENB), and

the third indicates the specimen number.

3. ^ Tests

To ensure that the weld variations represented only those due to the weave procedure, we measured both the tensile strength (at room temperature)

and CVN toughness (at -51 °C) of the undiluted weld metal near the center of

the four welds. The results were compared with the mechanical property

requirements of MIL-E-23765/2C, confirming that the welding procedure produced high quality welds.

For macroscopic examination of the fusion, specimens from each weld were

cut transverse to the welding direction (fig. 4). The samples were polished

and macroetched with a 10 percent nital solution. However with this standard

technique, only a section of the fusion line could be examined.

Another technique was developed to reveal the fusion-line surface in

three dimensions: the four welds (described in Section II) made with the

type 100S-1 electrode and HY-80 base material, were duplicated using type 309

stainless steel electrode and HY-80 base material with the welding conditions

in tables 5 and 6. These welds were cut from the plate at a point just

outside the HAZ and were dipped in a solution consisting of 100 m& H

2 0

2 , 100

mS, H

2 0, and 14 mS, HF. This solution dissolved the remaining HY-80 plate and

HAZ region but did not dissolve the stainless steel weld metal. Since only

the weld metal was left, the fusion-line surfaces were completely exposed

(fig. 5).

Using these two macroscopic techniques, we were able to study the entire

contact surface between the weld metal and the base material.

Elastic-plastic fracture toughness tests were performed at room temperature on specimens from each of the four welds. Although the base material

thickness was 25 mm, the fracture toughness specimen thickness was reduced to

22 mm in the machining operation and to 15 mm when side grooved. For the

fusion-line specimens, the initial notch and fatigue precrack were located as closely as possible to the fusion line. Figure 6 shows two examples of notch

location with respect to the fusion line. For specimen 1 -CT-2 (fig. 6a), the

6
notch is on the weld metal side of the fusion line; for specimen 2-CT-1

(fig. 6b), the notch is on the base material HAZ side of the fusion line. It was extremely difficult to locate the tip of the fatigue precrack precisely

at the fusion line. However, if the fusion line is truly the weakest region

of the weld, then the precrack does not have to be exactly on the fusion line

for the crack to grow along the fusion line.

For comparison, the fracture toughness of the base material and center

of the weld were measured. In the one base material specimen, the notch was

in the T-L plane, that is, parallel to the direction of welding. For the two

weld metal specimens, the notch was in the center of the weld metal parallel

to the direction of welding.

Both SENB and CT specimen configurations were employed in the toughness

testing (see fig. 7). Some of the CT specimens were side grooved 20 percent

(10 percent on each side) in the hope that the side grooves would limit the

crack growth to the region of the fusion line. To produce a fatigue precrack

with a relatively straight crack front, the sides of the SENB specimens (not

side grooved) were compressed a total of 2 percent (1 percent on each side)

[14].

Two examples of the fatigue precracks produced by the different test

procedures and specimen configurations are shown in figure 8. Figure 8a

shows a compact tension specimen that was side grooved prior to fatigue precracking. The fatigue precrack front is very irregular and does not meet

ASTM E8 1 3 requirements. Figure 8b shows an SENB specimen that yielded

locally in compression prior to fatigue precracking. The fatigue crack front

is relatively straight and is acceptable, according to ASTM E8l 3 . Note also

that the welding flaw had no apparent effect on the growth of the fatigue

precrack, but it did have a decided effect on the growth of the overload

crack, increasing crack growth in the vicinity of the flaw.

The tests were run on a servohydraulically controlled test machine with

a load capacity of 100 kN. A computer-aided interactive test procedure was used that calculated the J-integral according to the ASTM E8l 3 test procedure. The change in crack length from the unloading compliance was determined

at frequent intervals during the test [15].

After the fracture toughness tests, the fracture surfaces were examined

with a scanning electron microscope (SEM) to determine where the crao:

7
initiated and how it propagated. In the examination, we assumed that the

overload crack initiated at the place where the stretch zone is a minimum

[

1

6 ] . For example, consider the fracture surface of the specimen in

figure 9. The smooth region near the top is the fatigue precrack, and the

rough surface near the bottom is the subsequent fracture. The darker region

on the fatigue crack surface adjacent to the subsequent fracture is the

stretch zone. Its minimum width, near the center of the figure, is the

location of the origin of the fracture.

The location of the crack initiation with respect to the fusion line was determined by the size of the dimples at the initiation site. Figures 10 and

11 show the relative size of the dimples in the base metal and weld metal

specimens, respectively. Because the inclusion distribution determines the

dimple size and spacing and the weld metal has more inclusions than the base

metal, the weld metal fracture surface has more and smaller dimples than the

base metal fracture surface. Thus, it was possible to identify the material

at the crack initiation.

The HY-80 weld cross section used for the macrographic study was also

surveyed for microhardness . The hardness was measured along a line normal to

the fusion line and extending from the undisturbed base material, through the

HAZ and fusion line, and into the weld. The individual measurements were

taken at small intervals to indicate the trends in regions where the hardness

varied over a short distance (such as the HAZ and fusion line); larger

measurement intervals were chosen in the other regions. The fusion line was chosen as the origin when comparing the hardness traverses, with measurements

in the weld being assigned negative distances and the measurements in the

base material being assigned positive distances from this point. These

measurements were performed to determine the relative hardness (and so the

relative strength and toughness) of the fracture path.

A. RESULTS AND DISCUSSION

The tensile test results for the weld metal and the specification

requirements (MIL-E-23765/2C) are listed in table 7. Both the ultimate tensile strength (UTS) and yield strength in the weld metal met the specified

values. In the specimens (weld 3 and A) where the most severe weave pattern
was used, we noticed higher yield strength and UTS than in the other samples.

The 8 percent higher strength is attributed to the effects of greater base

material dilution with these procedures.

Table 8 presents the substantial compositional differences between the

base material and the welding electrode (taken from tables 3 and 4). When

the weld, already rich in manganese, is enriched with carbon and chromium

through dilution, the weld's carbon equivalent is increased substantially,

leading to a higher strength deposit.

The measured CVN test results for the weld metal, as well as the specification requirements (MIL-E-23765 ) , are shown in table 9. The weld metal

toughness values meet the specification requirements. The welds that had the

highest penetration into the side walls (welds 3 and 4) had 20 percent lower

absorbed energy than the other specimens. The same behavior can be seen in

the lateral expansion results. Again, the increased carbon equivalent due to

dilution by the base material is the cause, and the higher strength microstructure has a lower toughness.

The low-magnification photographs of transverse sections through the

four welds in figure 12 show the effects of horizontal arc oscillation on the

macrostructure . When oscillation was not used (fig. 12a), the beads were more rounded than when oscillation was used (fig. 12b, c,d). Figure 12 also

reveals that the HAZ thickness and fusion-line shape are the same for all

four welds. Metallographic studies did not show evidence of porosity or incomplete fusion in the weld.

Figure 13 shows the fusion-line surfaces for the four welds produced

using stainless steel filler material in order to characterize the texture of

the fusion line. This technique shows the entire surface clearly; only a line of which is revealed in conventional transverse sections. The stainless

steel weld metal has a different composition and therefore a different

viscosity and wetting character than the ferritic welds; however, the fusion

surface is representative of the welding arc action (which was duplicated)

and the weave pattern. The smooth fusion-line surface of weld 1 (stringer

bead technique) is vastly different than the convoluted fusion-line surface

for the three welds with the weave techniques. Only small differences are apparent among the three weave patterns; therefore, variable penetration into

joint walls and the difference between semiautomatic and adaptive control had

9
little effect on the fusion-line surfaces for these tests. The spacing and

height of the convolutions (ripples) should have a substantial effect on the

toughness of welds in which the fusion line is the critical crack path. Such

ripples tend to lengthen the fracture path (and thus increase the fracture

toughness) and tend to turn the running crack into the tougher weld or base

material

.

Careful examination of the fusion-line surfaces reveals small, flat areas that could be interpreted as insufficient fusion flaws. These regions,

comprising less than 2 percent of the fusion-line surface area, were not

observed on the weld radiographs of transverse sections. They could be the

weld flaws that were seen on several of the fusion-line fracture surfaces.

The fracture toughness test results are summarized in tables 10 and 11.

The values are referred to as J Q and K

Q (J), rather than J Ic and KIc , because

the results do not meet all the validity requirements specified in ASTM E8l 3

In most of the tests, the deviations from the standard test procedure are considered to have only minor effects on the results, and the values reported

should be representative of the true fracture toughness. However, the speci- mens with the highest Jq values, above 385 kJ/m2

, did not meet the minimum

thickness requirements and the values are dependent on specimen size.

Therefore, they are labeled invalid in the comments column. Although the

values are, strictly speaking, invalid, they still give an indication of high

toughness

.

For the specimens with the notch in the vicinity of the fusion line,

Kq(J) values varied from 1 80 to 315 MPaVm. The large scatter in the data

depended more on the exact location of the notch with respect to welding

flaws and the fusion line rather than on the weld weave procedure.

The specimens from the center of the welds (1-CT-WM and 4-CT-WM),

one from a straight weld and the second from a weave weld, were both very

tough, but the results were invalid for the specimen sizes used. The same result was found for the base metal specimen 4-CT-BM.

The fracture initiation location with respect to the fusion line was determined by studying the fracture surfaces, which are well defined in

figures 10 and 11. From this SEM examination, we can also follow the crack

propagation. The results are summarized in tables 1 0 and 11. Figure 14

shows one typical pattern, where the crack initiated in the base metal HAZ

10
and propagated into the weld metal (identified in table 10 by HAZ + WM)

.

This type of crack development occurred in specimens 2-CT-1 and 3 -CT-4.

Figure 15 shows another typical pattern, where the crack initiated at a weld

flaw and propagated in the weld (WM flaw). This type of pattern occurred in

specimens 1-CT-2, 1-CT-3, 3-CT-1, 3-CT-2, 1-SENB-1 and 3-SENB-1. Figure 16

shows the third typical pattern where the crack initiated and propagated in

the weld (WM). This pattern of crack development was present in specimens

H-CT-1 , 4-CT-2, 2-SENB-1, and 2-SENB-2.

The quantity of Jq is a measure of fracture initiation resistance, the

beginning of ductile tearing. Some of the specimens had low values of Jq,

but typically that low value can be attributed to the presence of welding

flaws. All the specimens exhibited an extremely steep tearing slope for

crack growth regardless of their Jq value. Figures 17, 18, and 19 are representative J-R curves of three different types of crack growth, distinct

from the three types of crack initiation described above. For specimen

1-CT-2, the crack initiated at a welding flaw and grew with an assist from

the flaw. As the crack front passed the flaw, it blunted. The J-R curve for

this specimen, figure 17, shows a relatively low value for Jq, a shallow

initial tearing slope followed by a much steeper tearing slope that corresponds to the crack blunting. For specimen 1-SENB-1, the crack initiated in

the weld near the fusion line at fine inclusions and grew through the weld.

The J-R curve, figure 18, shows a relatively good Jq value with a steep tearing

slope. For specimen 2-SENB-2, the crack initiated near the fusion line of

the weld, again at fine inclusions, and then grew through the weld metal. The

J-R curve, figure 19, shows a very high toughness--enough to invalidate the

test results for this size specimen. These test results indicate that

welding flaws near the fusion line do not promote brittle crack growth

behavior in HY-80 welds tested at room temperature.

Table 12 lists the hardness traverses across the four welds. The

traverse started in the weld (distance from the fusion line was chosed to

be negative) and continued through the HAZ into the base material. The

fusion line and HAZ immediately adjacent have hardnesses intermediate between

those of the weld and the rest of the HAZ. Therefore, the strength and

toughness values are also be expected to be intermediate, and this region is
not expected to be a critical failure path. These results support the

detailed fracture mechanics testing. 5.

CONCLUSIONS

1 . Four welds, produced with both semiautomat ically and adaptively guided

welding systems and several weave procedures, all met the mechanical

property requirements of MIL-E-23765/2C . The welding system itself did

not substantially affect the weld strength or toughness.

2. The most substantial differences among the welds were attributed to weld

penetration: the higher penetration level increased the weld strength

from 630 to 680 MPa and reduced the CVN toughness from 135 to 110 J.

3. Variations in crack initiation location and crack growth (J-R curve) were not related to the weld weave procedure. Study of other weld

parameters is needed to find the source of these variations.

4. The region near the fusion line of HY-80 welds has a lower toughness at room temperature than the HAZ or center of the weld, due to dilution by

the base material. However, the failure mechanism is ductile in all cases with reasonably good values for Jq, the energy to initiate a crack

.

5. The lowest values of Jq measured were associated with welding flaws.

6. The slopes of the tearing lines in the fracture- toughness tests were

quite steep, indicating a high resistance to crack growth.

7. Microhardness traverses indicated that the fusion line had hardness

values intermediate between those of the weld and the HAZ.

8. The entire fusion-line surface can be revealed by simulating the weld

with a stainless steel electrode and then dissolving the base material.

6.

ACKNOWLEDGMENTS

This effort was supported and monitored by Mr. C. A. Zanis of the D. W.

Taylor Naval Ship R&D Center under Program Element 62761 N, Task Area

SF 61541001 . The contributions of Mr. D. P. Vigliotti as technician on this

project are gratefully acknowledged. The base material was furnished by

Mr. R. Brooks of Phoenix Steel.

12
7. REFERENCES

[ 1 ] Military Specification, Electrodes and Rods - Welding, Bare, Solid,

or Alloyed Core, Low Alloy Steel , MIL-E-23765/2C (SH), June 1983.

Washington, D.C.: NAVSEA, Dept, of Navy.

[2] Military Specification, Electrodes, Welding, Mineral Covered, IronPowder, Low-Hydrogen, Medium, High Tensile and Higher-Strength Low

Alloy Steels , MIL-E-22200/ 1 OA (SH ) , July 1983. Washington D.C.: NAVSEA,

Dept, of Navy.

[3] Telford, R. T. 1965. New applications for gas metal arc welding and

short circuiting transfer, Welding Journal 49 : 169.

[4] Flax, R. W. ; Keith, R. E.; and Randall, M. D. 1971. Welding the HY Steels , ASTM STP 494. Philadelphia: American Society for Testing and Materials.

[5] Airco Shielded Metal Arc Electrode Data , Airco, 1985. Cleveland, Ohio.

[6] Atom Arc Iron Powder Low Hydrogen Electrodes ( AR-AA2-20M-579F ) , Alloy

Rods, Hanover, Pennsylvania.

[7] Hobart Electrodes, Wire - Flux (EW-518), Hobard Brothers Company, 1983.

Troy, Ohio.

[8] Weldirectory : of Electrodes for Carbon and Alloy Steel (M210), The

Lincoln Electric Company, 1985. Cleveland, Ohio.

[ 9 ] McKay Selector: A guide to the Selection and Use of McKay

Welding Electrodes and Wires (75F-2), Teledyne McKay, 1984. York,

Pennsylvania.

[10] Military Specification, Steel Plate, Alloy, Structural, High Yield

Strength (HY-80 and Hy-100) , MIL-S-1 621 6J (SH) , April 1981. Washington,

D.C. NAVSEA, Dept, of Navy.

[11] Certified Material Test Report from Phoenix Steel Corporation, Claymont,

Delaware, for Heat 54994 HY-80 base material used in this study.

[12] Savage, W. F.; Nippes E. F.; and Szekeres, E. S. 1976. A study of weld

interface phenomena in a low alloy steel. Welding Journal 55: 260-S.

[13] Baeslack, W. A., Ill; Lippold, J. C.; and Savage, W. F. 1979. Unmixed

zone formation in austenitic stainless steel weldments. Welding Journal

58: 168-S.

13
[14] Towers, 0. L. and Dawes, M. G. 1985. Welding Institute research on the

fatigue precracking of fracture toughness specimens. Elastic-Plastic

Fracture Test Methods: User's Experience . ASTM STP 856, eds. E. T.

Wessel and F. J. Loss: 23. Philadelphia: American Society for Testing

and Materials.

[15] Tobler, R. L . ; Read, D. T.; and Reed, R. P. 1981. Strength/toughness

relationship for interstitially strengthed AISI 304 stainless steels at

4 K temperature. Fracture Mechanics: Thirteenth Conference . ASTM STP

743, ed. R. Roberts: 250. Philadelphia: American Society for Testing

and Materials.

[16] Kobayashi, H.; Nakamura, H.; and Nakazawa, H. 1983. Comparison of J lc

test methods recommended by ASTM and JSME. Elastic-Plastic Fracture

Test Methods: User’s Experience . ASTM STP 856, eds. E. T. Wessel and F.

J. Loss: 3-22. American Society for Testing and Materials.
15

Figure

1.

Weld

joint

configuration

and

dimensions.
Figure 2. Plate alignment on the welding table.

16
Figure 3- View from the end of the weld illustrating

the torch to promote good fusion with both

the 20° rotation of

sides of the joint.

17
CD g

CO

<D

LLI _l<o

05Oz coo

03 Q

03

T

18

Figure

4.

Length

and

notch

orientation

of

specimens

removed

from

the

four

welds
igure 5. View of the stainless steel weld metal showing the fusionline surface after the HY-80 base plate was dissolved.

19
Figure 6. Two examples of notch location, deviating slightly from the fusion-line.

20
COMPACT TENSION SPECIMEN

3 POINT BEND SPECIMEN

Figure 7. Sing le-edge-notch bend and compact tension specimen configurations.

21
Weld

Raw

22

Figure

8.

Changes

in

fatigue

crack

shape

produced

by

precompression.
J Test

Fracture

Stretch

Zone

Fatigue

Precrack

t Weld Flaw 0.5 mm

Figure 9. Fracture surfaces showing crack initiation.

23
Figure 10. Fracture surface dimple size in the base metal.

24
£Gi

F igure 1 1 . Fracture surface dimple size in the weld.

25
Weld 3 Weld 4

Figure 12. Macrostructures of cross sections of the four welds.

26
27

Figure

13.

View

of

the

fusion-line

surface.
Large Dimples

(HAZ) J TestFracture

Stretch Zone

Fine Dimples

( WM) J Test

Fracture

50 f±m

Figure 14. Macrograph showing a fracture pattern of initiation

in the HAZ and propagation in the weld.

28
Fatigue

Precrack

Stretch

Zone

J Test

Fracture

0.5 mm

Figure 15. Macrograph showing a fracture pattern of initiation

at a weld flaw and propagation in the weld.

29
20 (JLm

Figure 16. Macrograph showing a fracture pattern of initiation

in the weld and propagation in the weld.

30
31
Figure 18. Representative J-R curve for crack initiation in the weld

32
33
Table 1 . Comparison of CVN absorbed energy for welds and base material

Mater ial Requ irement

Toughness

Achieved

GMA electrode

SMA electrode

48 J at -51 °C

48 J at -51 °C

81 -153 J at -51 °C

48-103 J at -51 °C

HY -80 base material 48 J at -84 °C 205 J at -84 °C

Table 2 Summary of the experimental design

Weld Torch Oscillation Penetration Welding System

Amplitude

1 No Low Semiautomatic

2 Yes Low Semiautomatic

3 Yes High Semiautomatic

4 Yes High Adap tive

34
Table 3- Chemical composition and mechanical properties of 25-mm-thick

HY-80 plate (MIL-S-1 621 6J ) used for these tests

Chemical Composition

Elements Weight %

c 0.14

Mn 0.28

Si 0.29

P 0.008

S 0.002

Cr 1 .32

Ni 2.42

Mo 0.30

A1 0.022

V 0.002

Ti 0.001

Cu 0.1 6

As 0.007

Sb 0.001

Sm 0.019

Mechanical Properties

Yield strength

Tensile strength

Elongation

Reduction of area

CVN*

CVN

625 MPa

740 MPa

48.5% in 50 mm

70 . 0 %

T - 268 J at - 1 8°C

B - 257 J at - 1 8°C

T - 204 J at -84 °C

B - 206 J at -84 °C

*CVN = Charpy V-notch

35
Table Chemical composition and mechanical properties of 1.2 mm diameter

MIL 100S-1 welding electrode*

Chemical Composition

Elements Weight %

c 0.068

Mn 1 .58

Si 0.34

P 0.008

S 0.003

Cr 0.1

1

Ni 1 .72

Mo 0.33

A l 0.01

V 0.01

Ti 0.01

Zr 0.02

Cu 0.03

Mechanical Properties

Yield strength 620 MPa

Tensile strength 740 MPa

Elongation 25.7% in 50 mm

Reduction of area 68.7 %

CVN T - 140 J at -1 8°C

B - 68 J at -51 °C

tested according to MIL-E-23765/2C

36
Table 5. Summary of constant welding conditions for the first and

subsequent passes for each of the test plates

Pass 1 ( root) 2-10

Filler Metal 100S - 1 1 oos-i

Shielding gas Ar-5 % C0 2 Ar-5 % COo

Torch Gap (mm) 22 22

Weave Width (mm) 10 10

Current (A) 220 280-300

Voltage (V) 25 28-30

Travel Speed (mm/min) 122 190

Heat input (kJ/mm) 2.7 2.6

37
Table 6. Typical weld weave widths for the various weld passes in welds 2 to 4

Weave Width

Pass 1 2 3 4 5 6 7 8 9

Osc

.

strok e ( mm) 9.4 10.1 9.4 9.4 9.4 9.4 10.4 10.4 1 1 .7

Osc

.

speed (mm/min) 550 710 710 710 710 710 1000 1000 1 1 40

38
Table 7. Tensile properties of the weld metal

Yield Strength Ultimate Tensile Elongation

Weld 0.2 % (MPa) Strength (MPa) (%)

1 (stringer) 628 707 30

2 (weave) 624 707 30

3 (weave) 694 746 27.5

4 (weave) 678 753 32

Specification 606-703 19

* (MIL-E -2 3765/2C)

minimum required

Table 8. Compositional differences between base material and welding

electrode

Base Weld ing

Element Material Electrode

(wt. %) (wt. %)

C 0.14 0.068

Mn 0.28 1 .58

Cr 1 .32 0.11

39
Table 9. Charpy V-notch data for the weld metal at -51 °C

Specimen Absorbed Lateral

Energy * Expansion

(J) ( mm)

1 a ( str inger) 121 1 .3

1 b ( stringer

)

118 1.3

1 c (str inger) 156 1.7

1 d ( stringer

)

155 1 .8

1 e ( str inger) 1 22 134 average 1 .4 1 .5 average

2 a (weave)

2 b (weave)

2 c (weave)

2 d (weave)

2 e (weave)

149

171

144

1 47

1 44

1 .8

1 .8

1 .7

1 .6

1 52 average 1 .

7

1 .7 average

3 a (weave) 110 1 .2

3 b (weave) 128 1 .4

3 c (weave) 125 1 .5

3 d (weave) 99 1 .0

3 e (weave) 1 1 7 1 1 6 average 1 .

3

1 .3 average

4 a (weave) . 98 1 .3

4 b (weave) 120 1.6

4 c (weave) 115 1 .5

4 d (weave) 90 1 .2

4 e (weave) 95 104 average 1 .3 1 .4 average

* Specification requirement: 47 J.

40
Table 10. Fracture toughness test results for 25-mm compact-tension

specimens

Spec imen Notch

Location

J

Q

( kJ/m^

)

Kq( J ) (MPa Vm)Initiation

Site

Comments

1-CT-1 FL 370 276 WM

1 -CT-2 FL 160 180 WM flaw

1-CT-3 FL 180 190 WM flaw

1 -CT-WM WM 600 352 WM invalid*

2-CT-1 FL 370 275 HAZ + WM

2-CT-2 FL 200 204 HAZ + WM

3-CT-1 FL 180 1 82 WM flaw

3-CT-2 FL 210 207 WM flaw

3-CT-3 FL 250 240 HAZ

3-CT-4 FL 230 219 HAZ + WM

4-CT-1 FL 1 70 188 WM

4-CT-2 FL 190 200 WM

4-CT-WM WM 400 289 WM invalid*

4-CT-BM BM 640 358 BM

% invalid

* Spec imen

ASTM E 8 1

3

thickness and the

requirements.

slope of the tearing line fail to meet

41
Table 1 1

.

spec imens

Fracture toughness test results for 25-mm single-edge-notch bend

Spec imen Notch

Location

J 0

( kJ/m^

)

Kq(J)

(MPa/m)

Initiation Comments

Site

1-SENB-1 FL 300 248 WM flaw

1 -SENB-2 FL 480 315 HAZ

2-SENB-1 FL 310 252 WM

2-SENB-2 FL 400 287 WM

3-SENB-1 FL 210 209 WM flaw

4-SENB-1 FL 230 219 FL flaw

4-SENB-1 FL 320 257 HAZ + WM Flaw

FL - Fusion-Line

42
Table 12. Microhardness traverse across the fusion line of the four welds

Distance from Hardness (Vickers-500

)

Fusion Line (mm) Weld 1 Weld 2 Weld 3 Weld 4

WM -1.5 232 263 252

-1 .0 287 315 366 275

-0.5 287 206

-0.4 202 21 4 263

-0.3 287 198 301 1 91 -0.2 281 339 315 21 4 -0.1 301 275 301 214

FL - 264 301 269 287

HAZ 0.1 366 348 315 315

0.2 348 366 301 275

0.3 263 206 301 275

0.4 287 241 348 263

0.5 275 366 331 275

1 .0 287 366 275 241

1 .5 263 385 366 21 4

BM 2.0 21 4 191 171 263

2.5 191 184

3.0 177 198 21 4 3.5 222 254 1 91 4.0 194

4.5 194 214

5.0 165 184

43
NBS-114A (REV. 2-80

U.S. DEPT. OP COMM. BIBLIOGRAPHIC DATA

SHEET (See i nstruction s)

. PUBLICATION OR

REPORT NO.

NBSIR 86-3043

2. Performing Organ. Report No Publication Date

April 1986

4. TITLE AND SUBTITLE

Fusion Line Shape versus Toughness in HY-80 GMA welds

5. AUTHOR(S)

T. A. Siewert, R. E. Trevisan, and P. T. Purtscher

6. PERFORMING ORGANIZATION (If joint or other than NBS, see in struction s) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State, ZIP)

D. W. Taylor Naval Ship R&D Center

Department of the Navy

10. SUPPLEMENTARY NOTES

J Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bi bliography or literature survey, mention it here) This report describes the effect of the electrode weave procedure on both the fusionline shape and toughness of reduced-gap gas metal arc (GMA) welds. To study the

variation in toughness, four GMA welds were produced in 25-mm-thick HY-80 plate

using MIL 100S-1 electrode and following the weld procedure listed in the electrode

specif ication, MIL-E-23765/2C. These four welds were used to compare stringer beads

with various weave procedures using both manually controlled and adaptively con- trolled welding systems. All-weld-metal Charpy V-notch (CVN) and tensile tests con- firmed that these welds met the requirements for MIL 100S-1 electrodes. An increase

in strength from 630 to 680 MPa and a decrease in toughness from 1 35 to 110 J were observed for the weave patterns with the greatest penetration into the base material.

Testing of compact tension and single-edge-notch bend specimens at room temperature

produced Kq(j) fusion line toughness values of 180 to 315 MPa- m, with little variation due to weave procedures. Microstructural examination of the crack- initiation

region indicated failures occurred and propagated in the HAZ or in the weld metal,

rather than along the fusion line. An etching technique revealed a small number of

planar flaws on the fusion-line surface that were not detected with radiography. We

conclude that weave procedure substantially modifies the fusion-line shape, can change the weld strength, and has little effect on fusion-line or weld toughness.

12. KEY WORDS (S/x to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) electrode weaving; fracture toughness ; fusion line; gas metal arc welding;

HY-80 ; mechanical properties; weld properties.

13.

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